Transcript of APPENDIX 1.A: ALLOY X DESCRIPTION
Attachment 5: HI-STORM 100 FSAR Proposed Changed Pages.HI-STORM 100
FSAR Proposed Rev. 134 REPORT HI-2002444 1.A-1
APPENDIX 1.A: ALLOY X DESCRIPTION 1.A ALLOY X DESCRIPTION 1.A.1
Alloy X Introduction Alloy X is used within this licensing
application to designate a group of stainless steel alloys. Alloy X
can be any one of the following alloys:
• Type 316 • Type 316LN • Type 304 • Type 304LN • Duplex Stainless
Alloy S32205 [1.A.3]
Qualification of structures made of Alloy X is accomplished by
using the least favorable mechanical and thermal properties of the
entire group for all MPC mechanical, structural, neutronic,
radiological, and thermal conditions. The Alloy X approach is
conservative because no matter which material is ultimately
utilized, the Alloy X approach guarantees that the performance of
the MPC will meet or exceed the analytical predictions. The duplex
stainless steel is deemed to be extremely resistant to stress
corrosion cracking (SCC) in marine environments. However, its
properties begin to deteriorate rapidly at temperatures above 600
oF reaching an uncertain state of fragility at 887 oF [1.A.4] and
above. Therefore, this material shall be used only if the metal
temperature of the MPC shell can be assured to remain below the
limit in Table 1.A.6 under all normal operating modes [1.A.3].
Likewise, under short term and accident conditions, such as the
“inlet duct blockage” scenario, the maximum metal temperature of
duplex stainless steel must be held below the limit in Table 1.A.6.
For other stainless steels listed as members of Alloy X above, the
design temperature limits in Table 2.2.3 remain unmodified. This
appendix defines the least favorable material properties of Alloy
X. 1.A.2 Alloy X Common Material Properties Several material
properties do not vary significantly from one Alloy X constituent
to the next. These common material properties are as follows:
• density • specific heat • Young's Modulus (Modulus of Elasticity)
• Poisson's Ratio
The values utilized for this licensing application are provided in
their appropriate chapters.
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1.A.3 Alloy X Least Favorable Material Properties The following
material properties vary between the Alloy X constituents:
• Design Stress Intensity (Sm) • Tensile (Ultimate) Strength (Su) •
Yield Strength (Sy) • Coefficient of Thermal Expansion (α) •
Coefficient of Thermal Conductivity (k)
Each of these material properties are provided in the ASME Code
Section II [1.A.1]. Tables 1.A.1 through 1.A.5 provide the ASME
Code values for each constituent of Alloy X along with the least
favorable value utilized in this licensing application. The ASME
Code only provides values to - 20oF. The design temperature of the
MPC is -40oF to 725oF as stated in Table 1.2.2. Most of the
above-mentioned properties become increasingly favorable as the
temperature drops. Conservatively, the values at the lowest design
temperature for the HI-STORM 100 System have been assumed to be
equal to the lowest value stated in the ASME Code. The lone
exception is the thermal conductivity. The thermal conductivity
decreases with the decreasing temperature. The thermal conductivity
value for -40oF is linearly extrapolated from the 70oF value using
the difference from 70oF to 100oF. The Alloy X material properties
are the minimum values of the group for the design stress
intensity, tensile strength, yield strength, and coefficient of
thermal conductivity. Using minimum values of design stress
intensity is conservative because lower design stress intensities
lead to lower allowables that are based on design stress intensity.
Similarly, using minimum values of tensile strength and yield
strength is conservative because lower values of tensile strength
and yield strength lead to lower allowables that are based on
tensile strength and yield strength. When compared to calculated
values, these lower allowables result in factors of safety that are
conservative for any of the constituent materials of Alloy X.
Further discussion of the justification for uUsing the minimum
values of coefficient of thermal conductivity has the effect of
reducing the heat rejection rate from the canister which is given
in Chapter 3conservative. The maximum and minimum values are used
for the coefficient of thermal expansion of Alloy X. The maximum
and minimum coefficients of thermal expansion are used as
appropriate in this submittal to support a conservative safety
evaluation. However, for any internal interference assessment the
actual values of coefficients of thermal expansion from the ASME
Code or Table 1.A.4 will be used. Figures 1.A.1-1.A.5 provide a
graphical representation of the varying material properties with
temperature for the Alloy X materials. 1.A.4 References [1.A.1]
ASME Boiler & Pressure Vessel Code Section II, 1995 ed. with
Addenda through 1997. [1.A.2] ASME Boiler & Pressure Vessel
Code Section II, 2013 ed. with Addenda through 2014 [1.A.3] ASME
Code Case N-741 (2013)
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HI-STORM 100 FSAR Proposed Rev. 134 REPORT HI-2002444 1.A-3
[1.A.4] C. Örnek, D. Engelberg, S. Lyon and T. Ladwein, "Effect of
“475°C Embrittlement” on the Corrosion Behaviour of Grade 2205
Duplex Stainless Steel Investigated Using Local Probing Techiques,"
Corrosion Management Magazine, no. 115, pp. 9-11, 2013..
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HI-STORM 100 FSAR Proposed Rev. 134 REPORT HI-2002444 1.A-4
Table 1.A.1 ALLOY X AND CONSTITUENT DESIGN STRESS INTENSITY (Sm)
vs. TEMPERATURE
Temp. (oF) Type 304 Type 304LN Type 316 Type 316LN
Duplex Stainless
Alloy X (minimum of constituent
values)
-40 20.0 20.0 20.0 20.0 31.7 20.0 100 20.0 20.0 20.0 20.0 31.7 20.0
200 20.0 20.0 20.0 20.0 31.7 20.0 300 20.0 20.0 20.0 20.0 30.6 20.0
400 18.7 18.7 19.3 18.9 29.4 18.7 500 17.5 17.5 18.0 17.5 28.7 17.5
600 16.4 16.4 17.0 16.5 28.4 16.4 650 16.2 16.2 16.7 16.0 28.3 16.0
700 16.0 16.0 16.3 15.6 - 15.6 750 15.6 15.6 16.1 15.2 - 15.2 800
15.2 15.2 15.9 14.9 - 14.9
Notes: 1. Source: Table 2A on pages 314, 318, 326, and 330 of
[1.A.1] for Type
316/316LN/304/304LN. 2. Units of design stress intensity values are
ksi. 3. Design stress intensity values have been derived based on
the basis established in
Mandatory Appendix 2 page 924 and 925 which essentially states that
the stress intensity value at temperature is the minimum of
one-third of the tensile strength or two-thirds of the yield
strength at temperature.
4. Maximum temperature of use for duplex stainless steel under both
long term storage and
short term / accident conditions is noted in Table 1.A.6.
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HI-STORM 100 FSAR Proposed Rev. 134 REPORT HI-2002444 1.A-5
Table 1.A.2 ALLOY X AND CONSTITUENT TENSILE STRENGTH (Su) vs.
TEMPERATURE
Temp. (oF) Type 304 Type 304LN Type 316 Type 316LN
Duplex Stainless
5]
values)
-40 75.0 (70.0) 75.0 (70.0) 75.0 (70.0) 75.0 (70.0) 95 (95) 75.0
(70.0)
100 75.0 (70.0) 75.0 (70.0) 75.0 (70.0) 75.0 (70.0) 95 (95) 75.0
(70.0)
200 71.0 (66.2) 71.0 (66.2) 75.0 (70.0) 75.0 (70.0) 95 (95) 71.0
(66.2) 300 66.0 (61.5) 66.0 (61.5) 73.4 (68.5) 70.9 (66.0) 91.7
(91.7) 66.0 (61.5) 400 64.4 (60.0) 64.4 (60.0) 71.8 (67.0) 67.1
(62.6) 88.2 (88.2) 64.4 (60.0) 500 63.5 (59.3) 63.5 (59.3) 71.8
(67.0) 64.6 (60.3) 86.1 (86.1) 63.5 (59.3) 600 63.5 (59.3) 63.5
(59.3) 71.8 (67.0) 63.1 (58.9) 85.2 (85.2) 63.1 (58.9) 650 63.5
(59.3) 63.5 (59.3) 71.8 (67.0) 62.8 (58.6) 85.0 (85.0) 62.8 (58.6)
700 63.5 (59.3) 63.5 (59.3) 71.8 (67.0) 62.5 (58.4) - 62.5 (58.4)
750 63.1 (58.9) 63.1 (58.9) 71.4 (66.5) 62.2 (58.1) - 62.2 (58.1)
800 62.7 (58.5) 62.7 (58.5) 70.9 (66.2) 61.7 (57.6) - 61.7 (57.6)
Notes: 1. Source: Table U on pages 437, 439, 441, and 443 of
[1.A.1] for Type
304/304LN/316/316LN. 2. Units of tensile strength are ksi. 3. The
ultimate stress of Alloy X is dependent on the product form of the
material (i.e.,
forging vs. plate). Values in parentheses are based on SA-336
forged materials (type F304, F304LN, F316, and F316LN) or SA-182
forged material (S32205), which are used solely for the one-piece
construction MPC lids. All other values correspond to SA-240 plate
material
4. Table U on page 529 of [1.A.2] for Duplex Stainless Steel S32205
5. Maximum temperature of use for duplex stainless steel under both
long term storage and
short term / accident conditions is noted in Table 1.A.6..
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Table 1.A.3
Temp. (oF) Type 304 Type 304LN Type 316 Type 316LN
Duplex Stainless
Alloy X (minimum of constituent
values)
-40 30.0 30.0 30.0 30.0 65.0 (70.0) 30.0 100 30.0 30.0 30.0 30.0
65.0 (65.1) 30.0 200 25.0 25.0 25.8 25.5 57.8 (62.2) 25.0 300 22.5
22.5 23.3 22.9 53.7 (57.9) 22.5 400 20.7 20.7 21.4 21.0 51.2 (55.2)
20.7 500 19.4 19.4 19.9 19.4 49.6 (53.4) 19.4 600 18.2 18.2 18.8
18.3 47.9 (51.6) 18.2 650 17.9 17.9 18.5 17.8 46.9 (50.5) 17.8 700
17.7 17.7 18.1 17.3 - 17.3 750 17.3 17.3 17.8 16.9 - 16.9 800 16.8
16.8 17.6 16.6 - 16.6 Notes: 1. Source: Table Y-1 on pages 518,
519, 522, 523, 530, 531, 534, and 535 of [1.A.1] for
Type 304/304LN/316/316LN. 2. Units of yield stress are ksi. 3.
Table Y-1 on page 676 and 677 of [1.A.2] for Duplex Stainless Steel
S32205. Values in
paratheses are based on SA-182 forged material (S32205) which is
used solely for the one-piece construction MPC lids. All other
values correspond to SA-240 plate material.
4. Maximum temperature of use for duplex stainless steel under both
long term storage and
short term / accident conditions is noted in Table 1.A.6.
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Table 1.A.4 ALLOY X AND CONSTITUENT COEFFICIENT OF THERMAL
EXPANSION vs. TEMPERATURE
Temp. (oF)
4]
Alloy X Maximum
Alloy X Minimum
-40 8.55 8.54 7.1 8.55 8.547.1 100 8.55 8.54 7.1 8.55 8.547.1 150
8.67 8.64 7.3 8.67 8.647.3 200 8.79 8.76 7.5 8.79 8.767.5 250 8.90
8.88 7.6 8.90 8.887.6 300 9.00 8.97 7.8 9.00 8.977.8 350 9.10 9.11
7.9 9.11 9.107.9 400 9.19 9.21 8.0 9.21 9.198.0 450 9.28 9.32 8.1
9.32 9.288.1 500 9.37 9.42 8.3 9.42 9.378.3 550 9.45 9.50 8.4 9.50
9.458.4 600 9.53 9.60 8.4 9.60 9.538.4 650 9.61 9.69 8.5 9.69
9.618.5 700 9.69 9.76 8.6 9.76 9.698.6 750 9.76 9.81 8.7 9.81
9.768.7 800 9.82 9.90 8.8 9.90 9.828.8
Notes: 1. Source: Table TE-1 on pages 590 and 591 of [1.A.1], for
Type 304/304LN/316/316LN. 2. Units of coefficient of thermal
expansion are in./in.-oF x 10-6. 3. Table TE-1 on page 753 of
[1.A.2] for Duplex Stainless Steel S32205. 4. Maximum temperature
of use for duplex stainless steel under both long term storage
and
short term / accident conditions is noted in Table 1.A.6.
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Table 1.A.5 ALLOY X AND CONSTITUENT THERMAL CONDUCTIVITY vs.
TEMPERATURE
Temp. (oF) Type 304
4]
values)
-40 8.23 6.96 7.83 6.96 70 8.6 7.7 8.2 7.7 100 8.7 7.9 8.3 7.9 150
9.0 8.2 8.6 8.2 200 9.3 8.4 8.8 8.4 250 9.6 8.7 9.1 8.7 300 9.8 9.0
9.3 9.0 350 10.1 9.2 9.5 9.2 400 10.4 9.5 9.8 9.5 450 10.6 9.8 10.0
9.8 500 10.9 10.0 10.2 10.0 550 11.1 10.3 10.5 10.3 600 11.3 10.5
10.7 10.5 650 11.6 10.7 10.9 10.7 700 11.8 11.0 11.2 11.0 750 12.0
11.2 11.4 11.2 800 12.2 11.5 11.6 11.5
Notes: 1. Source: Table TCD on page 606 of [1.A.1] for Type
304/304LN/316/316LN.
2. Units of thermal conductivity are Btu/hr-ft-oF.
3. Table TCD on page 773 of [1.A.2] for Duplex Stainless Steel 4.
Maximum temperature of use for duplex stainless steel under both
long term storage and
short term / accident conditions is noted in Table 1.A.6.
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Table 1.A.6 DUPLEX STAINLESS STEEL TEMPERATURE LIMITS†
Parameter Value Long Term, Normal Condition Design Temperature
Limits (Long-Term Events) (o F)
525
Short-Term Events, Off-Normal, and Accident Condition Temperature
Limits (o F)
650
† These temperature limits take precedence over those in Table
2.2.3
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FIGURES 1.A.1 through 1.A.5 INTENTIONALLY DELETED
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Type Criteria Basis FSAR Reference
Minimum Cooling Time: 3 years (Intact ZR Clad Fuel) 8 years (Intact
SS Clad Fuel)
- Section 2.19
Max. Fuel Assembly Weight: (including non-fuel hardware and DFC, as
applicable)
1,720 lb. for fuel assemblies that do not require fuel
spacers,
otherwise 1,680 lb.
176.8 in. - Section 2.1.9
8.54 in. - Section 2.1.9
BWR Fuel Assemblies: Type Various - Sections 2.1.9 and 6.2 Max.
Burnup 65,000 MWD/MTU - Section 2.1.9 Max. Enrichment Varies by
fuel design - Section 2.1.9, Table 2.1.4
Max. Decay Heat/ MPC†.
- Section 4.4
Minimum Cooling Time: 3 2 years (Intact ZR Clad Fuel) 8 years
(Intact SS Clad Fuel)
Section 2.1.9
830 lb. (intact fuel) - Section 2.1.9
Max. Fuel Assembly Length (Unirradiated Nominal)
176.5in. - Section 2.1.9
5.85 in. - Section 2.1.9
Handling Loads 115% of Dead Weight CMAA #70 Section 2.2.1.2
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Proposed Rev. 143 REPORT HI-2002444 2.III-2
non-mechanistic). The basis for the lateral deflection limit in the
active fuel region, θ, is provided in [2.III.6.1] as Error! Objects
cannot be created from editing field codes. where δ is defined as
the maximum total deflection sustained by the basket panels under
the loading event and w is the nominal inside (width) dimension of
the storage cell. The limiting value of θ is provided in Table
2.III.4. The above deflection-based criterion has been used
previously in the HI- STAR 180 Transportation Package [2.III.6.2]
to qualify similar Metamic-HT fuel baskets. ii. Thermal The design
and operation of the HI-STORM 100 System with the MPC-68M must meet
the intent of the review guidance contained in ISG-11, Revision 3
[2.0.8] as described in Subsection 2.0.1. All applicable material
design temperature limits in Section 2.2 and 4.3 continue to apply
to the MPC-68M. Temperature limits of MPC-68M fuel basket and
basket shim materials are specified in Table 4.III.2. The MPC-68M
is designed for both uniform and regionalized fuel loading
strategies as described in Subsection 2.0.1. The regions for the
MPC-68M are given in Table 2.III.1. Additionally, ta
quarter-summetric heat load pattern has been defined for MPC-68M as
shown in Figure 2.III.1. The same temperature limits apply to this
configuration. iii. Shielding Same as Subsection 2.0.1. iv.
Criticality Same as Subsection 2.0.1 with the clarifications
herein. Criticality control is maintained by the geometric spacing
of the fuel assemblies and spatially distributed B-10 isotope in
the Metamic-HT. No soluble boron is required in the MPC-68M water.
The minimum specified boron concentration in the Metamic-HT
purchasing specification must be met in every lot of the material
manufactured. No credit is taken for burnup. Enrichment limits are
delineated in Table 2.III.2. v. Confinement Same as Subsection
2.0.1 vi. Operations
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1 0.5*
2 0.5*
3 0.5*
4 0.5
5 1.2
6 1.2
7 0.5
8 0.5*
9 0.5*
10 0.5
11 1.2
12 0.4
13 0.4
14 1.2
15 0.5
16 0.5*
17 0.5
18 1.2
19 0.4
20 0.4
21 0.4
22 0.4
23 1.2
24 0.5
25 0.5*
26 1.2
27 0.4
28 0.4
29 0.4
30 0.4
31 0.4
32 0.4
33 1.2
34 0.5*
35 0.5*
36 1.2
37 0.4
38 0.4
39 0.4
40 0.4
41 0.4
42 0.4
43 1.2
44 0.5*
45 0.5
46 1.2
47 0.4
48 0.4
49 0.4
50 0.4
51 1.2
52 0.5
53 0.5*
54 0.5
55 1.2
56 0.4
57 0.4
58 1.2
59 0.5
60 0.5*
61 0.5*
62 0.5
63 1.2
64 1.2
65 0.5
66 0.5*
(kW)
* Note: This figure provides per cell allowable heat loads for
MPC-68M with all UNDAMAGED FUEL assemblies. For MPC-68M with
DAMAGED FUEL and/or FUEL DEBRIS stored in this location (in a DFC),
the per cell allowable heat load of the cell is limited to 0.35
kW.
Figure 2.III.1 Per Cell Allowable Heat Loads (kW) for
Quarter-Symmetric Pattern -
MPC-68M
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free from condensation and gross environmental contaminants. The
cleanliness requirements and inspections during fabrication and
fuel loading operations also ensure that the MPC has minimal
surface debris and impurities. Tests on Metamic-HT Extensive tests
[1.III] have been conducted to establish material properties of
Metamic-HT including its corrosion-resistance characteristics. The
Metamic-HT specimens were used for corrosion testing in
demineralized water and in 2000 ppm boric acid solution. The tests
concluded that the Metamic-HT panels will sustain no discernible
degradation due to corrosion when subjected to the severe thermal
and aqueous environment that exists around a fuel basket during
fuel loading or unloading conditions. Aluminum Alloy Aluminum alloy
used in the fuel basket shims are hard anodized to achieve the
desired emissivity specified in Supplement 43.III. The anodizing is
an electrolytic passivation process used to increase the thickness
of the natural oxide layer on the surface of metal parts. Anodizing
increases corrosion resistance and wear resistance of the material
surface. There is no mechanistic process for the basket shims with
hard anodized surface to react with borated water or demineralized
water during fuel loading operation. Under the long-term storage
condition, the basket shims are exposed to dry and inert helium
with no potential for reaction. Finally, to ensure safe fuel
loading operation, the operating procedure described in Chapter 8
provides for the monitoring of hydrogen gas in the area around the
MPC lid prior to and during welding or cutting activities. Although
the aluminum surfaces (Metamic-HT fuel basket and aluminum basket
shims) are anodized, there is still a potential for generation of
hydrogen in minute amounts when immersed in spent fuel pool water
for an extended period. Accordingly, as a defense-in-depth measure,
the lid welding procedure requires purging the space below the MPC
lid prior to and during welding or cutting operation to eliminate
any potential for formation of any combustible mixture of hydrogen
and oxygen. Following the completion of the MPC lid welding and
hydrostatic testing, the MPC-68M is drained and dried. After the
completion of the drying operation, there is no credible mechanism
for any combustible gases to be generated within the MPC-68M.
3.III.4.2 Positive Closure Same as in Subsection 3.4.2. 3.III.4.3
Lifting DevicesAttachment Points The structural analyses of the
lifting devices attachment points in Subsection 3.4.3 (including
all paragraphs) are bounding for the MPC-68M for the following
reasons:
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During a non-mechanistic tip-over event, the fuel assemblies exert
a lateral force on the fuel basket panels as the overpack impacts
the ground and decelerates. The lateral force causes the fuel
basket panels to deflect potentially affecting the spacing between
stored fuel assemblies. To maintain the fuel in a subcritical
configuration, a deflection limit for the fuel basket panels is set
in Subsection 2.III.0.1, which is supported by the criticality
safety analysis in Supplement 6.III. Here a finite element analysis
is performed using ANSYS to demonstrate that the maximum lateral
deflection in the fuel basket panels under a bounding deceleration
of 60g is less than the limit specified in Section 2.III.0.1. The
60g input deceleration is bounding because it exceeds the design
basis deceleration limit of 45g for the non-mechanistic tip over of
the HI-STORM storage overpack (see Subsection 3.III.4.10) and for
the horizontal drop of the HI-TRAC transfer cask (see Subsection
3.4.9), and it matches the design basis lateral deceleration limit
of 60g for the HI-STAR transport cask [1.1.3] for future
considerations. The analysis methodology presented in this
subsection is identical to the methodology used in [2.III.6.2] to
qualify the F-37 fuel basket. As shown in Figure 3.III.1, a
representative slice of the MPC-68M fuel basket, consisting of a
smaller end section and a full section, is modeled in detail
including the contained fuel assemblies and supporting basket
shims. The fuel basket panels are modeled with SOLSH190 solid shell
elements. The basket shims and each fuel assembly are modeled with
SOLID45 solid elements. The mass density assigned to the fuel
assemblies corresponds to the maximum BWR fuel assembly weight per
Table 2.1.22, except at the 16 cell locations along the basket
perimeter where Damaged Fuel Containers are permitted. At these 16
locations, the mass density corresponds to the maximum weight of a
BWR fuel assembly plus DFC per Table 2.1.22. Standard contact pairs
using CONTA173/TARGE170 elements are defined at the interfaces of
fuel assembly/basket panel, shim/basket panel, and between stacked
basket panels including all the intersecting slot locations. At the
perimeter corners, the intersecting basket panels are bonded
together in the finite element model, and the strength properties
of the corner most elements are then adjusted depending on whether
there is a full length weld at that location. At corner locations
that are not welded full length (see licensing drawing in Section
1.5), the elastic modulus of the corner elements is reduced to 1%
of the MGV in Table 1.III.2 to effectively eliminate the joint’s
shear and moment carrying capacity. The fuel basket material model
is implemented with true stress-true strain multi-linear isotropic
hardening plasticity model. An elastic material model is used for
the basket shims since no plastic deformation is expected. To
accommodate large plastic deformation in the fuel basket panels,
sufficiently small element sizes (< 0.40 in) are used and 9
integration points through the thickness are specified. A
sensitivity study was performed in [2.III.6.2] to confirm that the
panel stresses and displacements obtained using solid shell
elements are converged and comparable to those obtained using 5
solid elements through the thickness of the panel. The 60g
deceleration is applied to the model with the basket in the
so-called 0º orientation (see Figure 3.III.5). This orientation is
chosen for analysis because it maximizes the lateral load on a
single basket panel, which in turn maximizes the lateral deflection
of the panel. In the 0º orientation, the amplified weight of each
stored fuel assembly (during the 60g impact event) bears entirely
on one basket panel. Conversely, in the 45º orientation, the
amplified weight of each stored fuel assembly is equally supported
by two basket panels. The difference in loading between these two
basket orientations is pictorially shown in Figure 3.III.5, where
“m” denotes
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based on Charpy V-notch absorbed energy (CVE) correlations for
steels. The estimated value is consistent with the range for
aluminum alloys, which is 20 to 50 MPa√m or 18.2 to 45 ksi√in per
Table 3 of [3.III.4]. Next the minimum crack size, amin, for crack
propagation to occur is calculated below using the formula for a
through-thickness edge crack given in [3.1.5]. Although the formula
is derived for a straight-edge specimen, the use of the peak
stress, σmax, at a notch in the fuel basket panel (instead of the
average stress in the panel as required by the formula) essentially
compensates for the geometric difference between the basket panel
and the specimen. Moreover, the maximum size of a pre-existing
crack (1/16”) in the fuel basket panel is less than 1/6th of the
basket panel thickness (0.40”). Thus, the assumption of a
through-thickness edge crack is very conservative. The result
is
=
1.12 2
= 1.3980.752
and the safety factor against crack propagation (based on a 1/16”
minimum detectable flaw size) is
=
= 22.412.0
The calculated minimum crack size is more than 12 22 times greater
than the maximum possible pre-existing crack size in the fuel
basket (based on 100% surface inspection of each panel). The large
safety factor ensures that crack propagation in the MPC-68M fuel
basket will not occur due to the non-mechanistic tipover event.
3.III.4.4.3.2 Elastic Stability and Yielding of the MPC-68M Fuel
Basket under
Compression Loads (Load Case F3 in Table 3.1.3) Under certain
conditions, the fuel basket plates may be under direct compressive
load. Although the finite element simulations can predict the onset
of an instability and post-instability behavior, the computation in
this subsection uses (the more conservative) classical instability
formulations to demonstrate that an elastic instability of the
basket plates is not credible. A solution for the stability of the
fuel basket plate is obtained using the classical formula for
buckling of a wide bar [3.III.1]. Material properties are selected
corresponding to a metal temperature of 325ºC375ºC, which bounds
the computed metal temperatures anywhere in the fuel basket (see
Table 4.III.3). The critical buckling stress for a pin-ended bar
is:
inksiKIC 30=
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where h is the plate thickness, a is the unsupported plate length,
E is the Young’s Modulus of Metamic-HT at 325ºC375ºC, ν is
Poisson’s Ratio (use 0.3 for this calculation) From the drawings in
Section 1.5, h = 0.40 in, a = 6.05 in, and E = 7,8506,125 ksi
(Table 1.III.2). Then, the classical critical buckling stress is
computed as 31.01424.199 ksi, which exceeds the yield strength of
the material (9.425 ksi) at 375ºC. This demonstrates that basket
plate instability by elastic buckling is not possible. 3.III.4.5
Cold Same as in Subsection 3.4.5. 3.III.4.6 HI-STORM 100 Kinematic
Stability under Flood Condition (Load Case A
in Table 3.1.1) The stability evaluation of the HI-STORM 100
overpack under flood conditions in Subsection 3.4.6 bounds the
scenario of a loaded MPC-68M inside a HI-STORM overpack. The
previous analysis is bounding because it uses as input the empty
weight of the HI-STORM overpack (i.e., no MPC inside) combined with
the maximum CG height from Table 3.2.3. 3.III.4.7 Seismic Event and
Explosion Since there are no physical changes to the HI-STORM
overpacks and the MPC-68M reduces the CG height of the loaded
HI-STORM overpacks, relative to those analyzed in Chapter 3, the
seismic event and explosion analyses presented in Subsection 3.4.7
(including all paragraphs) bound the scenario of a loaded MPC-68M
inside a HI-STORM overpack. 3.III.4.8 Tornado Wind and Missile
Impact (Load Case B in Table 3.1.1 and Load
Case 04 in Table 3.1.5) The results for the post-impact response of
the HI-STORM 100 overpack in Subsection 3.4.8 for the combination
of tornado missile plus either steady tornado wind or instantaneous
tornado pressure drop bound the results for a loaded MPC-68M inside
a HI-STORM overpack. The results are bounding because they are
calculated assuming a lower bound weight for the loaded HI-STORM
and an upper bound CG height (as compared to a loaded MPC-68M
inside a HI- STORM). In addition, since the MPC-68M does not
require any physical changes to the HI-STORM overpacks or the
HI-TRAC transfer casks for MPC loading, the missile penetration
analyses presented in Subsection 3.4.8 remain valid.
( ) ( ) 2
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Same as in Subsection 3.4.12 and with the following supplementary
information provided herein. 3.III.4.12.1 Metamic-HT Considerations
Metamic-HT has been extensively tested as indicated in Supplement
1.III. Testing has included extensive tests for creep, irradiation
and corrosion to ensure long-term fuel basket performance under
normal conditions of storage. The Metamic-HT is also not
susceptible to structural fatigue and brittle fracture under long
term conditions of storage. Corrosion is discussed further in
Subsection 3.III.4.1. Creep and boron depletion are further
discussed below. i) Fuel Basket Creep The Metamic sourcebook
contains data on the testing to determine the creep characteristics
of the Metamic-HT under both unirradiated and irradiated
conditions. A creep equation to estimate a bounding estimate of
total creep as a function of stress and temperature is also
provided. The creep equation developed from this test provides a
conservative prediction of accumulated creep strain by direct
comparison to measured creep in unirradiated and irradiated
coupons. The creep equation for Metamic-HT that bounds all measured
data (tests run for 20,000 hours) is of the classical exponential
form in stress and temperature (see Supplement 1.III)[1.III.3],
which is written symbolically as ε = f(σ,T). Creep in the MPC-68M
fuel basket will not be a reactivity modifier because the basket is
arrayed in the vertical orientation. The lateral loading of the
fuel basket walls is insignificant and hence no mechanistic means
for the basket panels to undergo lateral deformation from creep
exists, even if the panel material were susceptible to creep. The
creep effect would tend to shorten the fuel basket under the
self-weight of the basket. An illustrative calculation of the
cumulative reduction of the basket length is presented below to
demonstrate the insignificant role of creep in the MPC-68M fuel
basket. The in-plane compressive stress, σ, at height x in the
basket panel is given by σ = ρ(H-x) (3.III.1) where: ρ = weight
density of Metamic-HT H = height of the fuel basket Using the above
stress equation, the total creep shrinkage, δ, is given by
(3.III.2)
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where: T = panel’s metal temperature (conservatively assumed to be
350375°C for a period of 60 years) H = height of the basket
(conservatively assumed to be 200 inches) Using the creep equation
(provided in Supplement 1.III[1.III.3]) and performing the above
integration numerically yields δ = 0.138095 inch. In other words,
the computed shrinkage of the basket is less than 0.06948% of its
original length. Therefore, it is concluded that for the vertical
storage configuration the creep effects of the MPC-68M fuel basket
are insignificant due to absence of any meaningful loads on the
panels. Therefore, creep in the Metamic-HT fuel basket is not a
matter of safety concern. ii) Fuel Basket Boron Depletion The
similarities between Metamic-HT and Metamic (classic) neutron
absorbers and their exposure to the same long-term conditions of
storage in the HI-STORM 100 system provide a logical basis to
expect negligible neutron absorber boron depletion in Metamic-HT.
However, to assure criticality safety during worst case design
basis conditions over the 40-year design life, the analysis
discussed in Subsection 6.III demonstrates that the boron depletion
in the Metamic- HT is negligible over a 50-year duration. Thus,
sufficient levels of boron are present in the fuel basket to
maintain criticality safety over the 40-year design life of the
MPC. 3.III.4.12.2 Basket Shim Considerations: i) Basket Shim Creep
Like the fuel basket, the basket shims are not subject to any
significant loading during storage. The ability of the basket shims
(made of a creep resistant aluminum alloy) has been evaluated and
qualified in Docket No. 71-9325 [2.III.6.2] for transport
applications where the stress level (in horizontal configuration)
is significant. Therefore, in light of the minuscule stress levels
from self-weight in long-term storage, creep is ruled out as a
viable concern for the basket shims. ii) Basket Shim Corrosion
Basket shim corrosion is discussed in Subsection 3.III.4.1.
3.III.4.13 Design and Service Life Same as in Subsection 3.4.13.
3.III.5 FUEL RODS Same as in Section 3.5. 3.III.6 SUPPLEMENTAL
DATA
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TABLE 3.III.4 MAXIMUM DISPLACEMENT IN MPC-68M FUEL BASKET
Maximum Lateral
(Note 1)
Maximum Allowable Value of θ (from Table 2.III.4) Safety
Factor
8.91.008 × 10-34 0.005 5.624.96 Notes: 1. See Subsection 2.III.0.1
for definition of θ.
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FIGURE 3.III.2: FIRST PRINCIPALTRUE STRESS DISTRIBUTION IN MPC-68M
FUEL BASKET UNDER 60G LOAD
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4.4.2 [deleted] 4.4.3 Test Model The HI-STORM thermal analysis is
performed on the FLUENT [4.1.2] Computational Fluid Dynamics (CFD)
program. To ensure a high degree of confidence in the HI-STORM
thermal evaluations, the FLUENT code is benchmarked using data from
tests conducted with casks loaded with irradiated SNF
([4.1.3],[4.1.7]). The benchmark work is archived in QA validated
Holtec reports ([4.1.5],[4.1.6]).These evaluations show that the
FLUENT solutions are conservative in all cases. In view of these
considerations, additional experimental verification of the thermal
design is not necessary. 4.4.4 Maximum and Minimum Temperatures
4.4.4.1 Maximum Temperatures The 3-D model from the previous
subsection is used to determine temperature distributions under
long-term normal storage conditions for an array of cases covering
PWR and BWR fuel storage in uniform and regionalized loading
configurations. For this purpose one bounding MPC design in each of
the two fuel classes – MPC-68 for BWR and MPC-32 for PWR – are
analyzed and results obtained and summarized in this subsection.
For a bounding evaluation the MPCs are assumed to be emplaced in a
limiting overpack (HI-STORM 100S Version B). The HI-STORM 100S
Version B is the limiting overpack by virtue of the inlet and
outlet vents design. Compared to two other overpack designs (i.e.,
HI-STORM 100 and HI-STORM 100S), the HI-STORM 100S Version B has
smaller inlet and outlet vents. Thus Version B vent airflow
resistances are bounding. Also, the HI-STORM 100S Version B is the
shortest of the overpacks. This reduces the chimney height which
minimizes the driving head for air flow. Because the HI- STORM 100S
Version B will have the least cooling air flow, it will yield
bounding results. A cross-reference of HI-STORM thermal analyses is
provided in Table 4.4.5. Under regionalized loading, an array of
runs covering a range of regionalized storage configurations
specified in Chapter 2 (X=0.5 to X=3) are analyzed. The results are
graphed in Figures 4.4.6 and 4.4.7 for PWR and BWR fuel storage
respectively. Based on this array of runs the fuel storage
condition corresponding to X = 0.5 is determined to be limiting for
both PWR and BWR MPCs. Accordingly HI-STORM MPC and overpack
temperatures are reported for this storage condition in Tables
4.4.6 and 4.4.7. Damaged fuel is canestirezed in damaged fuel
containers (DFCs) before long-term storage. Each MPC type has
designated locations for placement of DFCs, as described in Chapter
2. Particularly, the DFCs are placed for storage in basket
peripheral locations. The presence of DFCs impedes helium flow
through it. However, since the DFCs are placed in the cold
peripheral locations, they do not control the peak cladding
temperature. Moreover, as a substantial fraction of basket cells
are occupied by intact fuel, the overall effect of DFC fuel storage
on the heat dissipation from the basket
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is small. To account for its impact, a heat load penalty described
in Tables 4.4.3 and 4.4.4 is placed on the DFC locations under both
uniform and regionalized heat load patterns. It should be noted
that the 3-D FLUENT cask model incorporates the effective
conductivity of the fuel assembly submodell. Therefore the FLUENT
models report the peak temperature in the fuel storage cells Thus,
as the fuel assembly models include the fuel pellets, the FLUENT
calculated peak temperatures are actually peak pellet centerline
temperatures which bound the peak cladding temperatures with a
margin. The following observations can be derived by inspecting the
temperature field obtained from the thermal models:
• The fuel cladding temperatures are below the regulatory limit
(ISG-11 [4.1.4]) under all storage scenarios (uniform and
regionalized) in all MPCs.
• The maximum temperature of the basket structural materials are
within their design limits.
• The maximum temperature of the neutron absorbers are below their
design limits.
• The maximum temperatures of the MPC pressure boundary materials
are below their design
limits.
• The maximum temperatures of concrete is within the guidance of
the governing ACI Code (see Table 4.3.1).
The above observations lead us to conclude that the temperature
field in the HI-STORM System with a loaded MPC containing heat
emitting SNF complies with all regulatory temperature limits. In
other words, the thermal environment in the HI-STORM System is in
compliance with Chapter 2 Design Criteria. 4.4.4.2 Minimum
Temperatures In Table 2.2.2 of this report, the minimum ambient
temperature condition for the HI-STORM storage overpack and MPC is
specified to be -40°F. If, conservatively, a zero decay heat load
with no solar input is applied to the stored fuel assemblies, then
every component of the system at steady state would be at a
temperature of -40°F. Low service temperature (-40oF) evaluation of
the HI- STORM is provided in Chapter 3. All HI-STORM storage
overpack and MPC materials of construction will satisfactorily
perform their intended function in the storage mode at this minimum
temperature condition. 4.4.4.3 Effects of Elevation The reduced
ambient pressure at site elevations significantly above the sea
level will act to reduce the ventilation air mass flow, resulting
in a net elevation of the peak cladding temperature. However,
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Table 4.4.3
MAXIMUM PERMISSIBLE HEAT LOAD[deleted] FOR LOCATIONS WITH DFCs
UNDER
UNIFORM LOADING
MPC Type Decay Heat for Locations with Damaged Fuel Assemblies and
Fuel Debris (kW)
MPC-24E/24EF 1.114
MPC-32/32F 0.718
MPC-68/68FF/68M 0.393
Table 4.4.4
MAXIMUM PERMISSIBLE HEAT LOAD FOR LOCATIONS WITH DFCs UNDER
REGIONALIZED LOADING
MPC Type Decay Heat for Locations with Damaged Fuel Assemblies and
Fuel Debris (kW)
MPC-24E/24EF 0.75*q2 (Note 1)
MPC-32/32F 0.65*q2 (Note 1)
MPC-68/68FF/68M 0.75*q2 (Note 1) Note 1: q2 is the maximum
permissible heat load allowed for intact fuel in Region 2.
[deleted]
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MW = minimum water flow rate (lb/hr) Cpw = water heat capacity
(Btu/lb-°F) Tmax = maximum MPC cavity water mass temperature (must
be less than 212oF) Tin = MPC water inlet temperature Q =
Coincident fuel decay heat in the canister (Btu/hr)
For example, the MPC cavity water temperature limited to 150°F, MPC
water inlet temperature at 125°F and design basis maximum heat load
(36.9 kW, approximately 125,908 Btu/hr), the water flow rate
computes as 5038 lb/hr (10.1 gpm). 4.5.3 MPC Temperatures During
Moisture Removal Operations 4.5.3.1 Vacuum Drying Operation The
initial loading of SNF in the MPC requires that the water within
the MPC be drained, fuel dried and the water replaced with helium.
Vacuum drying of fuel is conducted by evacuating the MPC after
completion of MPC draining operation. For MPCs containing Moderate
Burnup Fuel (MBF) assemblies only, this operation may be carried
out using the vacuum drying method up to the threshold heat loads
defined in Table 4.5.1. In this Table threshold heat loads Q1 and
Q2 are defined wherein Q1 is the threshold heat load for vacuum
drying operations without time limits and Q2 is the threshold heat
load for time-limited vacuum drying. The requirements and limits
for moisture removal are provided in LCO 3.1.1 of the HI-STORM 100
CoC and are specific to the amendment to which the HI-STORM 100
System is being loaded. To minimize fuel temperatures during vacuum
drying operations the HI-TRAC annulus must be water filled. At heat
loads greater than the threshold heat load defined above, the peak
cladding temperature cannot be maintained below the ISG-11,
Revision 3 limit of 570oC for MBF under a vacuum condition of
infinite duration. Under this scenario, cycles of vacuum drying
resulting in heatup followed with cooling by helium are performed
until drying criteria is achieved. Similarly, vVacuum drying of
MPCs containing one or more High Burnup Fuel (HBF) assemblies is
also not permitted under time limits. The peak cladding temperature
for drying HBF must be maintained below ISG-11 Rev 3 limit of
400oC. It must be noted that the permissible time for
heatup/cooldown cycles is a function of canister specific heat
loads. At lower heat loads the duration of vacuum drying cycles is
higher. The thermal model defined below must be used for
heatup/cooldown cycles for site-specific canister heat load maps.
It must be ensured per ISG-11 Rev 3 that the repeated thermal
cycling is limited to less than 10 cycles, with cladding
temperature variations less than 65oC (117oF) each cycle. High
burnup fFuel drying must bemay also be conducted by using forced
helium drying (FHD) process as discussed in Section 4.5.3.2. A 3-D
FLUENT thermal model of the MPC is constructed in the same manner
as described in Section 4.4. The principal input to this model is
the effective conductivity of fuel under vacuum drying operations.
To reasonably bound vacuum drying operations the effective
conductivity of fuel is computed assuming the MPC is filled with
water vapor at a very low pressure (1 torr) for the
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entire duration of vacuum drying17. The methodology for computing
the effective conductivity is given in Section 4.4.1. To ensure a
conservative evaluation the thermal model is incorporated with the
following assumptions:
i. Threshold heat load Q1, defined in Table 4.5.1, is assumed and
steady-state condition reached under Q1 results in vacuum drying of
only MBF without time limits.
ii. Threshold heat load Q2, defined in Table 4.5.1, is assumed and
a transient calculation is performed to determine the permissible
vacuum drying time under Q2. The transient calculation is started
assuming the MPC has reached 212oF boiling temperature in the
operational step preceding vacuum drying (i.e. water blow down
operations). The vacuum drying clock starts when the MPC is
drained.
iii. The external surface of the MPC shell is postulated to vary
linearly from 100oC (212oF) normal boiling temperature of water at
the top to 111oC (231oF) elevated pressure boiling temperature at
the bottom to account for the hydrostatic head.
iv.iii. The bottom surface of the MPC is insulated. v.iv. MPC
internal convection heat transfer is suppressed. vi.v. Top surface
of the MPC is in communicative contact with air (Table 2.2.2).
Natural
convection and radiation cooling from the MPC top is included in
the thermal model. The principle objective of the vacuum drying
analysis is to ensure that fuel temperatures are below ISG-11, Rev.
3 temperature limits (See Table 4.3.1). Under threshold heat load
Q1 the results and margins are tabulated in Table 4.5.5. Under the
time limited threshold heat load Q2 the peak cladding temperature
plot is shown in Figure 4.5.2. The results under the scenarios Q1
and Q2 (with appropriate time limit) show that ISG-11, Rev. 3
limits are met with ample margins. 4.5.3.2 Forced Helium
Dehydration To dry the MPC cavity using a Forced Helium Dehydration
(FHD) system, a conventional, closed loop dehumidification system
consisting of a condenser, a demoisturizer, a compressor, and a
pre- heater is utilized to extract moisture from the MPC cavity
through repeated displacement of its contained helium, accompanied
by vigorous flow turbulation. A vapor pressure of 3 torr or less is
assured by verifying that the helium temperature exiting the
demoisturizer is maintained at or below the psychrometric threshold
of 21oF for a minimum of 30 minutes. See Appendix 2.B for detailed
discussion of the design criteria and operation of the FHD system.
The FHD system provides concurrent fuel cooling during the moisture
removal process through forced convective heat transfer. The
attendant forced convection-aided heat transfer occurring during
operation of the FHD system ensures that the fuel cladding
temperature will remain below the applicable peak cladding
temperature limit for normal conditions of storage, which is well
below the high burnup cladding temperature limit 752oF (400oC) for
all combinations of SNF type, burnup,
17 This is conservative as the MPC pressure is progressively
lowered below ambient pressure to facilitate moisture
removal. Near the end of the vacuum drying operation the pressure
is substantially lowered to approximately 1 torr to facilitate the
30-minute 3-torr vacuum rebound test followed by backfilling of the
MPC with helium.
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Table 4.5.1
Drying Method Fuel Burnup Threshold Heat LoadNote 1 Time
Limits
Vacuum Drying MBF Q1 None
Vacuum Drying MBF MPC Heat Load > Q1 and ≤ Q2
Yes (Note 2) (40 hrs)
Vacuum Drying HBF 36.9 kW Yes (Note 2)
FHD MBF and/or HBF 36.9 kW None Note 1: Threshold heat load iss are
defined below:as Q1 = 26 kW (Uniform) Note 2: Vacuum drying of the
MPC must be performed using cycles of the drying system, according
to the guidance contained in ISG-11 Revision 3 and as described in
Paragraph 4.5.3.1. The time limit for these cycles shall be
determined based on site specific conditionsQ2 = 30 kW
(Uniform).
Table 4.5.2
Component Weight (lbs) Heat Capacity
(Btu/lb-°F) Thermal Inertia
(Btu/°F) Water Jacket 7,000 1.0 7,000 Lead 52,000 0.031 1,612
Carbon Steel 40,000 0.1 4,000 Alloy-X MPC (empty)
39,000 0.12 4,680
Fuel 40,000 0.056 2,240 MPC Cavity Water* 6,500 1.0 6,500
26,032 (Total) * Conservative lower bound water mass.
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Table 4.5.4
Condition* Fuel in MPC Heat Load Reduction Factor * SCS
Required
1 All MBF 100% NO
2 One or More HBF ≤ 90% NO
3 One or More HBF > 90% YES
* The threshold heat load is obtained by multiplying the design
basis heat load per storage cell defined in Subsection 2.1.9.1 by
the reduction factor listed in this table.
Table 4.5.5
Threshold Heat LoadNote 1 Time Limit Temperature (oF)
Temperature
LimitNote 2 Margin (oF)
Q2 40 hrs 1035 1058 23 Notes:
1) Threshold heat load is defined in Table 4.5.1. 2) Temperature
limit of moderate burnup fuel shown. Vacuum drying of high burn-up
fuel
is not permitteddiscussed in (See Subsection 4.5.3.).
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Figure 4.5.2: PEAK CLADDING TEMPERATURE CURVE UNDER VACUUM
DRYING
OPERATIONS AT THRESHOLD HEAT LOAD Q2INTENTIONALLY DELETED
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SUPPLEMENT 4.III1
THERMAL EVALUATION OF THE MPC-68M
4.III.0 OVERVIEW The MPC-68M is a 68 cell BWR canister engineered
with a high B10 containing Metamic-HT basket for enhanced
criticality control. The MPC-68M is evaluated for storage in the
aboveground family of HI-STORM overpacks. For a bounding evaluation
an MPC-68M emplaced in the most flow resistive HI-STORM 100S
Version B overpack2 is analyzed under normal, off-normal and
accident conditions. The evaluations described herein parallel
those of the aboveground HI-STORM cask contained in the main body
of Chapter 4 of this FSAR. In addition, a new heat load layout is
added which is referred to as the "Quarter Symmetric Heat Load"
(QSHL) pattern. In this pattern, the maximum permissible heat load
in each storage cell, q, is specific to its location within the
quadrant and is limited to a unique prescribed value given in
Figure 2.III.1. This QSHL pattern seeks to minimize the large
temperature differences between cladding temperatures in proximate
fuel assemblies and is especially suited for canisterizing of fuel
with widely varying specific heat loads such as at a plant
undergoing decommissioning. It should be noted that the QSHL
pattern is a special case of regionalized loading, but is
identified simply as “QSHL” to avoid confusion. To ensure
readability, the section in the main body of the chapter to which
each section in this supplement corresponds is clearly identified.
All tables in this supplement are labeled sequentially. 4.III.1
INTRODUCTION The information presented in this supplement is
intended to serve as a complement to the information provided in
the main body of Chapter 4. Except for the fuel basket and basket
support materials, the information in Chapter 4 that remains
applicable to the MPC-68M analysis is not repeated herein.
Specifically the following information in the main body of Chapter
4 is not repeated:
1. The thermal properties of materials in Section 4.2 applicable to
the MPC-68M. 2. The specifications for components in Section 4.3
applicable to the MPC-68M. 3. The descriptions of the thermal
modeling of the MPC and its internals, including fuel
assemblies, in Section 4.4 which are applicable in their entirety
to the MPC-68M.
4. The descriptions of the short-term loading operations, carried
out using the HI-TRAC transfer cask, in Section 4.5 applicable to
the MPC-68M.
1 For ease of supplement review the sections are numbered in
parallel with the main Chapter 4. 2 This approach is identical to
the HI-STORM thermal analysis in Section 4.4.
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As confirmed by appropriate supporting analyses, the heat rejection
capacity of the MPC-68M1 is equal to or better than its
counterparts (strictly speaking, much better because of the highly
conducting Metamic-HT fuel basket). This renders its resistance to
accident events such as fire with greater margins of safety.
4.III.2 THERMAL PROPERTIES OF MATERIALS2 The material properties
compiled in Section 4.2 of the FSAR provide the required
information, except for the material properties of Metamic-HT fuel
basket, aluminum basket shims3 and solid shims. The Metamic-HT and
shims thermo-physical properties data is provided in Table 4.III.1.
4.III.3 SPECIFICATIONS FOR COMPONENTS4 All applicable material
temperature limits in Section 4.3 of the FSAR continue to apply to
the MPC- 68M. Temperature limits of MPC-68M fuel basket and basket
shim materials is specified in Table 4.III.2. 4.III.4 THERMAL
EVALUATION FOR NORMAL CONDITIONS OF STORAGE5 4.III.4.1 Thermal
Model The MPC-68M thermal design is same as that of the currently
licensed MPC-68. It features a 68 cells capacity fuel basket for
storing BWR fuel. The basket is engineered with a bottom plenum by
providing flow holes, a top open plenum by providing an engineered
clearance and a peripheral downcomer to facilitate heat dissipation
by thermosiphon action. The MPC-68M is helium pressurized to same
backfill specifications as discussed below:
• Initial helium backfill pressure is defined in Chapter 4, Table
4.4.12 under uniform and regionalized loading based on
regionalization parameter X (Section 2.1.9), and.
• Initial helium backfill pressure is defined in Table 1.III.1
under QSHL pattern defined in Supplement 2.III.
The principal differences are in the basket material of
construction (Metamic-HT), the installation of aluminum basket
shims in the basket peripheral spaces and replacement of the cell
walls sandwich
1 Heat rejection capacity is defined as the amount of heat the
storage system containing an MPC loaded with CSF stored
in uniform storage will reject with the ambient environment at the
normal temperature and the peak fuel cladding temperature at
400ºC.
2 This section supplements Section 4.2. 3 The terms basket shims
and extruded shims are interchangeably used in this chapter. 4 This
section supplements Section 4.3. 5 This section supplements Section
4.4.
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be supported on a subgrade at 77ºF. This is the same boundary
condition applied to the bottom of the ISFSI pad for the HI-STORM
100 modeling in Section 4.4. 4.III.4.2 Thermal Analysis The MPC-68M
has been designed to permit storage under the array of uniform and
regionalized heat loads defined in Chapter 2 as a function of the
regionalization parameter X. As shown in Chapter 4 the highest
cladding temperatures are reached under regionalized storage at X =
0.5. This scenario is co-incident with the maximum permissible MPC
heat load and therefore temperatures of other sub- systems (such as
fuel basket, MPC shell and overpack) also reach their highest
values. The fuel cladding temperature under long term storage in
HI-STORM is presented in Table 4.III.3.a The QSHL pattern has also
been analyzed using the same FLUENT model previously used in this
FSAR: no changes were made to the existing thermal model. The
selected heat loads in Figure 2.III.1 are suitably limited to
ensure that the peak cladding temperature in the MPC remains below
that in the governing MPC analyzed in this FSAR (MPC-32) under all
thermal scenarios. Thus the peak cladding temperature for the QSHL
pattern is limited by a previously analyzed and licensed MPC. Other
important safety aspects of the QSHL pattern are:
1. The hottest fuel assemblies are located in-board of the
peripheral locations in the basket so that the colder fuel in the
peripheral cells helps block the radiation emitted by the hottest
fuel assemblies.
2. The cell specific heat load, q, provided in Figure 2.III.1 is
the maximum value permitted for that location. In virtually every
case, the actual heat load in every cell will be lower than the
allowed limit, thus resulting in a lower cladding temperature field
overall than that computed herein.
3. The fuel cladding temperature for QSHL pattern under long term
storage in HI-STORM is
presented in Table 4.III.3.a. The predicted PCT is higher than that
for the scenario with decay heat based on regionalized parameter X
defined in Chapter 2. For this reason, QSHL pattern is adopted as
the licensing basis pattern for MPC-68M.
4. The PCT and basket temperatures under the QSHL pattern is lower
than that in the thermally governing case (MPC-32).
This QSHL scenario is adopted for demonstration of compliance with
the temperature and pressure limits set forth in this Supplement
and Chapter 2. The limiting scenario is analyzed and maximum
temperatures and pressures under normal storage tabulated in Tables
4.III.3b and 4.III.4. The results are below the Chapter 2 and
Supplement 4.III normal temperature and pressure limits. In
accordance with NUREG-1536 MPC-68M pressures are computed assuming
1% (normal), 10% (off-normal) and 100% (accident) rod ruptures with
100% rods fill gases and fission gases release in accordance
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with NUREG-1536 release fractions. The pressures are computed and
tabulated in Table 4III.4. The 100% rods rupture pressure is below
the accident design pressure (Table 2.2.1). 4.III.4.3 Engineered
Clearances to Eliminate Thermal Interferences To minimize thermal
stresses in load bearing members, the MPC-68M is engineered with
adequate gaps to permit free thermal expansion of the fuel basket
and MPC in axial and radial directions. In this subsection,
differential thermal expansions are evaluated to ensure the
adequacy of engineered gaps. The following gaps are
evaluated:
e.a. Fuel Basket-to-MPC Radial Gap f.b. Fuel Basket-to-MPC Axial
Gap g.c. MPC-to-Overpack Radial Gap h.d.MPC-to-Overpack Axial
Gap
The FLUENT thermal model articulated above provides the temperature
field in the HI-STORM overpack and MPC-68M from which the changes
in the above gaps are directly computed. The nominal cold gaps are
presented on the drawings in Section 1.5 and the corresponding
differential expansions under normal storage conditions are
presented in Table 4.III.8. The calculations show significant
margins against restraint to free-end expansion are available in
the design. 4.III.4.4 Evaluation of Fuel Debris Storage Fuel debris
is permitted for storage in up to eight peripheral cells under the
permitted uniform loading heat load limits specified in Section 2.4
of the Technical Specificationsheat load pattern shown in Figure
2.III.2. Although fuel debris is not required to meet cladding
temperature limits, its effect on fuel stored in the interior cells
must be assessed. Fuel debris in the canister is thermally
conservatively evaluated assuming a bounding debris configuration
and design heat load in all storage cells. The following
assumptions are adopted to maximize the computed cladding
temperatures:
1. The fuel debris is assumed to be completely pulverized and
compacted into a square prismatic bar enclosed by the damaged fuel
canister (DFC) with open helium space above it. In this manner the
height of the prismatic bar emitting heat is minimized resulting in
the maximization of lineal thermal loading (kw/ft) of the DFC and
co- incident local heating of the fuel basket and neighboring
storage cells.
2. Fuel debris assumed to be completely composed of UO2. As UO2 has
a lower conductivity relative to cladding, heat dissipation is
understated.
3. The fuel debris is assumed to block through flow of helium
inside the DFC. 4. All 16 peripheral storage locations (not just
the 8 permitted by CoC) are assumed to
contain fuel debris emitting maximum heat permitted by Technical
Specifications (CoC Appendix B, Section 2.4, Table 2.4-1Figure
2.4-2) and all interior cells are emitting design basis heat under
the uniform applicable heat loading storage scenario.
5. The MPC operating pressure is understated to minimize internal
convection heat transfer
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The results of the analysis are tabulated in Table 4.III.11. The
results support the following conclusions:
- Cladding temperature is substantially below the ISG-11, Rev. 3
limit. - MPC basket is below the design limit (Table 4.III.2) by
large margin. - MPC shell and Overpack metal temperatures are below
design limits (Table 2.2.3). - Overpack body and lid concrete are
well below design limits (Table 4.3.1).
4.III.5 THERMAL EVALUATION OF SHORT TERM OPERATIONS 4.III.5.1
HI-TRAC Thermal Model The HI-TRAC thermal model presented in
Section 4.5 is adopted for the evaluation of MPC-68M under short
term operations. 4.III.5.2 Maximum Time Limit During Wet Transfer
Operations As the MPC thermal inertia credited in the time-to-boil
calculations is bounded by the MPC-68M thermal inertia the
evaluation of wet transfer operations in Section 4.5 remains
applicable to the MPC-68M. Time-to-boil is calculated using the
same methodology described in Section 4.5.2. Table 4.III.13
summarizes the thermal inertia of the constituent components in the
loaded HI-TRAC transfer cask. Using the methodology described in
Section 4.5.2, the time-to-boil is provided at representative
initial temperatures for maximum QSHL in Table 4.III.14. This is an
example calculation for the maximum design basis heat load. The
same methodology can be adopted to determine the time-to-boil for
canisters loaded at lower heat loads. An alternate method using the
FLUENT thermal model described in Section 4.III.5.1 can be adopted
to evaluate the time for water within the MPC to boil using for
site-specific conditions. 4.III.5.3 MPC Temperature During Moisture
Removal Operations 4.III.5.3.1 Vacuum Drying Prior to helium
backfill the MPC-68M must be drained of water and demoisturized. At
the start of draining operation, both the HI-TRAC annulus and the
MPC are full of water. The presence of water in the MPC ensures
that the fuel cladding temperatures are lower than design basis
limits by large margins. As the heat generating region is uncovered
during the draining operation, the fuel and basket mass will
undergo a monotonic heat up from the initially cold conditions when
the heated surfaces were submerged under water. To limit fuel
temperatures demoisturization of the MPC-68M by the vacuum drying
method is permitted provided the HI-TRAC annulus remains water
filled during vacuum drying operations. To support vacuum drying
operations two limiting scenarios are defined below:
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Scenario A: The MPC-68M is loaded with Moderate Burnup Fuel
assemblies generating heat at the maximum permissible rate defined
in Chapter 2 under the bounding regionalized storage scenario X =
0.5.
Scenario B: The MPC-68M is loaded with one or more High Burnup Fuel
assemblies and the
MPC-68M decay heat is less than a conservatively defined threshold
heat load Q = 29 kW1.
To evaluate the above scenarios the vacuum drying analysis
methodology presented in Section 4.5 is adopted and an MPC-68M
specific thermal model constructed. The principal features of the
thermal model are as follows:
i. A bounding steady-state analysis is performed under the heat
loads defined in the scenarios above.
ii. The water in the HI-TRAC annulus is conservatively assumed to
be boiling under the hydrostatic head of water at the annulus
bottom (232oF).
iii. The bottom surface of the MPC is insulated. The thermal model
articulated above is used to compute the maximum cladding
temperature under the vacuum drying scenarios defined above. The
results tabulated in Table 4.III.5 are in compliance with the
ISG-11 temperature limits of Moderate Burnup Fuel (Scenario A) and
High Burnup Fuel (Scenario B). At heat loads greater than threshold
heat load defined above, the peak cladding temperature cannot be
maintained below the ISG-11, Revision 3 limit of 400oC for HBF
under a vacuum condition of infinite duration. Under this scenario,
cycles of vacuum drying resulting in heatup followed with cooling
by helium are performed until drying criteria is achieved. The
thermal model described above is used for heatup/cooldown cycles
for site-specific canister heat load maps. It must be ensured per
ISG-11 Rev 3 that the repeated thermal cycling is limited to less
than 10 cycles, with cladding temperature variations less than 65oC
(117oF) each. 4.III.5.3.2 Forced Helium Dehydration Evaluation of
Forced Helium Dehydration in Section 4.5 is applicable to MPC-68M.
4.III.5.3.3 Open Loop Low Pressure Drying (LPD) Method Dehydration
of the MPC cavity after draining off its bulk water can be carried
out using the vacuum drying method. Calculations show that below a
threshold heat load QL (Scenario B in Paragraph 4.III.5.3.1), the
peak cladding temperature of the used fuel is assured to remain
below the ISG-11 Rev 3 limit of 400oC for HBF regardless of the
duration of near vacuum conditions.
1 Threshold heat load is defined as the product of maximum loaded
assembly heat load rmax and the number of fuel storage cells
(n=68). Under this stipulation rmax must not exceed 0.426 kW.
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Therefore, at or below QL, the vacuum drying operation can be
continued for as long as necessary to achieve the target vacuum
pressure of 3 torr (or dew point of the contained helium gas <21
deg. F). At heat loads above QL, the vacuum drying operation cannot
be continued indefinitely lest the PCT rise above the ISG-11 Rev 3
limit. The duration t for which the state of near vacuum can be
held to keep the PCT below the ISG-11 Rev 3 limit is computed by a
transient analysis for the site-specific heat load using the FLUENT
model described in Paragraph 4.III.5.3.1. Accordingly, the vacuum
drying process is stopped at time t and the next phase, termed
"feed and bleed" is initiated. The feed and bleed process entails
introducing helium in sufficient quantity through the canisters’s
top vent opening to the cask cavity to raise the pressure to
approximately 500 mm Hg thus lowering the canister spatial
temperature by a modest amount, After that, a slow introduction of
helium through the top vent is begun along with a concomitant slow
withdrawal through the drain port. A nominal rate of helium "feed
and bleed” is provided in Table 4.III.10 . The slow bleed and feed
process is intended to create quiescent conditions in the cask
cavity which would facilitate the settling of the water vapor near
the bottom of the cask and stratification of the helium mass above
it. The bleeding of the gas from the bottom thus steadily dew
scavenges the water vapor from the cask reducing the helium mass’
relative humidity. The “dew point” of the exiting vapor (target ≤
21 deg. F) provides the definitive proof as to whether the canister
has been dried to the requisite level. The steady state peak
cladding temperature assuming helium at atmospheric pressure under
the most limiting pattern i.e. QSHL pattern, is presented in Table
4.III.10. The PCT during feed and bleed process is below the ISG-11
Rev 3 temperature limit of 400oC. The above process has the
principal benefit of avoiding cyclic heating and quenching of the
fuel associated with vacuum drying & flooding which is limited
in the number of permitted cycles and the related cyclic
temperature range by regulatory guidance to avoid high thermal
stresses. 4.III.5.4 Cask Cooldown and Reflood During Fuel Unloading
Operations Evaluation of cask cooldown and reflood operation in
Section 4.5 is applicable to MPC-68M. 4.III.5.5 HI-TRAC Onsite
Transfer Operation A 3D FLUENT thermal model of an MPC-68M emplaced
in a HI-TRAC transfer cask is constructed to evaluate the thermal
state of fuel under onsite transport in the vertical orientation1.
A bounding analysis is performed under the following conditions:
(i) Steady state maximum temperatures have reached.
(ii) The MPC-68M is loaded with fuel generating heat at the maximum
permissible level under the limiting regionalized storage scenario
X = 0.5Quarter Symmetric Heat Load (QSHL) pattern.
1 In accordance with Section 4.5 onsite transfer in the horizontal
orientation is not permitted.
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(iii) The HI-TRAC annulus is air filled. The scenario defined above
represents upper bound temperatures reached in the HI-TRAC without
the aid of any auxiliary cooling such as the Supplemental Cooling
System (SCS) defined in Section 4.5. The maximum cladding
temperatures computed using the thermal model articulated above are
tabulated in Table 4.III.6. As the cladding temperatures are below
the limiting High Burnup Fuel temperature limits mandated by ISG-11
[4.1.4] SCS cooling is not necessary for ensuring cladding safety
under onsite transfer operations involving for the MPC-68M
canister. Accordingly SCS cooling is not mandated in the MPC-68M
Technical Specifications. Additionally, the peak fuel cladding
temperatures are bounded by MPC-32 (Section 4.5). 4.III.5.6
Sensitivity Study In lieu of anodization of the extruded shims used
in the MPC-68M, they are passivated in water to form a thin oxide
layer. The emissivity of extruded shim surfaces is therefore
reduced and requires a thermal evaluation. In addition to the
above, the radial gap between the basket, extruded shims and MPC
shell is controlled using thin solid aluminum plates, which may be
inserted between the basket and extruded shims to meet the gap
criterion specified on the drawing (see Section 1.III.5). If the
gap criterion on the drawing is met, solid shims are not required.
These solid thin shim plates are made of aluminum and are supported
by the extruded shims. A thermal analysis is performed in this
subsection to determine the effect of these thin solid shim plates
and low emissive extruded shims. The following changes are made to
the thermal model discussed in previous sub-sections to study the
impact of the above mentioned design enhancements:
1. The panel notch gap on each side of the intersecting basket
panels is increased to 0.8mm. 2. The gap between the basket and
extruded shims is modeled with an effective thermal
conductivity. The effective thermal conductivity of the gap between
the basket and extruded shims is calculated based on a
two-dimensional CFD model. This 2-D model includes the solid shim
placed between the basket and extruded shims. A schematic of the
model is shown in Figure 4.III.1.
3. A conservatively lowerbound emissivity of 0.03 is used for the
passivated extruded shim surfaces.
4. Emissivity of solid shims is shown in Table 4.III.1. The solid
shims are conservatively modeled to be equidistant from the basket
wall and extruded shim wall. The effective thermal conductivity of
the gap between the basket and extruded shims with the presence of
solid shim plate bounds the scenario without the presence of solid
shims. The sensitivity study documented herein therefore considers
only the scenario with solid shim plate placed in the gap between
the basket and extruded shim. A sensitivity study is performed to
evaluate the most limiting thermal scenario with least margins to
fuel cladding temperature limit i.e. vacuum drying Scenario B
defined in Section 4.III.5.3. The results of the sensitivity study
to evaluate the effect of design enhancements made to the MPC
and
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its contents are reported in Table 4.III.12. The results
demonstrate that fuel temperature is well below its temperature
limit and is also bounded by the results based on the licensing
basis thermal model in Table 4.III.5. Therefore, the design
enhancements discussed in this subsection are bounded by the
licensing basis thermal analysis documented in this chapter. No
additional thermal analysis for other conditions (also considering
the large temperature margins to limits) is therefore warranted.
4.III.6 THERMAL EVALUATION OF OFF-NORMAL AND ACCIDENT CONDITIONS1
4.III.6.1 Off-Normal Conditions (a) Elevated Ambient Air
Temperature This off-normal event is defined in Paragraph 4.6.1.2.
The principal effect of elevated ambient temperature is a rise of
the HI-STORM 100 temperatures from the baseline normal storage
temperatures by the difference between elevated ambient and normal
ambient temperatures. The results of this event (maximum
temperatures and pressures) are provided in Table 4.III.15. The
results are below the off-normal condition temperature and pressure
limits (Tables 2.2.1, 4.III.2 and 2.2.3). As the normal storage
temperatures under MPC-68M storage in the HI-STORM 100 overpack are
bounded by the HI-STORM 100 System temperatures reported in Section
4.4, the temperatures under this event are likewise bounded by the
off-normal ambient evaluation in Section 4.6. (b) Partial Blockage
of Air Inlets This off-normal event is defined in Paragraph
4.6.1.3. The principal effect of partial inlet svent blockage is a
rise in the HI-STORM 100 annulus temperature rise in HI-STORM 100
System components from the baseline normal storage temperatures and
to leading order a similar rise in the MPC temperatures. As the
normal storage temperatures under MPC-68M storage in the HI-STORM
100 overpack are bounded by the HI-STORM 100 System temperatures
reported in Section 4.4, the temperatures under this event are
likewise bounded by the partial ducts blockage evaluation in
Section 4.6.Reasonably bounding evaluations in Paragraph 4.6.1.3
yield a certain rise in fuel cladding and component temperatures
due to this off-normal event. This temperature adder is applied to
HI-STORM MPC-68M storage temperature field in Table 4.III.3b and
presented in Table 4.III.15. The results yield substantial margins
for assuring safe storage of spent nuclear fuel, fuel basket and
MPC confinement boundaries. (c) Off-Normal Pressure 1 This section
supplements Section 4.6.
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This event is defined as a combination of (a) maximum helium
backfill pressure (Table 4.4.121.III.1), (b) 10% fuel rods rupture,
and (c) limiting fuel storage configuration. The principal
objective of the analysis is to demonstrate that the MPC off-normal
design pressure (Table 2.2.1) is not exceeded. The MPC-68M
off-normal pressure is reported in Table 4.III.4. The result1 is
below the off-normal design pressure (Table 2.2.1). 4.III.6.2
Accident Conditions (a) Fire Although the probability of a fire
accident affecting a HI-STORM 100 System during storage operations
is low due to the lack of combustible materials at an ISFSI, a
conservative fire event has been assumed and analyzed. The only
credible concern is a fire from an on-site transport vehicle fuel
tank. Under a postulated fuel tank fire, the outer layers of
HI-TRAC or HI-STORM overpacks are heated for the duration of fire
by the incident thermal radiation and forced convection heat
fluxes. The amount of fuel in the on-site transporter is limited to
a volume of 50 gallons. (i) HI-STORM Fire2 The fuel tank fire is
conservatively assumed to surround the HI-STORM Overpack.
Accordingly, all exposed overpack surfaces are heated by radiation
and convection heat transfer from the fire. Based on NUREG-1536 and
10 CFR 71 guidelines [4.III.2], the following fire parameters are
assumed:
1. The average emissivity coefficient must be at least 0.9. During
the entire duration of the fire, the painted outer surfaces of the
overpack are assumed to remain intact, with an emissivity of 0.85.
It is conservative to assume that the flame emissivity is 1.0, the
limiting maximum value corresponding to a perfect blackbody
emitter. With a flame emissivity conservatively assumed to be 1.0
and a painted surface emissivity of 0.85, the effective emissivity
coefficient is 0.85. Because the minimum required value of 0.9 is
greater than the actual value of 0.85, use of an average emissivity
coefficient of 0.9 is conservative.
2. The average flame temperature must be at least 1475°F (800°C).
Open pool fires typically
involve the entrainment of large amounts of air, resulting in lower
average flame temperatures. Additionally, the same temperature is
applied to all exposed cask surfaces, which is very conservative
considering the size of the HI-STORM cask. It is therefore
conservative to use the 1475°F (800oC) temperature.
3. The fuel source must extend horizontally at least 1 m (40 in),
but may not extend more than
3 m (10 ft), beyond the external surface of the cask. Use of the
minimum ring width of 1 meter yields a deeper pool for a fixed
quantity of combustible fuel, thereby conservatively maximizing the
fire duration.
1 Pressures relative to 1 atm absolute pressure (i.e. gauge
pressures) are reported throughout this section. 2 The HI-STORM
fire accident methodology is same as the generic methodology in
Section 4.6 of the HI-STORM 100 FSAR.
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The time constant of the cask body (i.e., the overpack) can be
determined using the formula:
Error! Objects cannot be created from editing field codes. where:
cp = Overpack Specific Heat Capacity (Btu/lb-oF) ρ = Overpack
Density (lb/ft3) Lc = Overpack Characteristic Length (ft) k =
Overpack Thermal Conductivity (Btu/ft-hr-oF) The concrete
contributes the majority of the overpack mass and volume, so we
will use the specific heat capacity (0.156 Btu/lb-oF), density (140
lb/ft3) and thermal conductivity (1.05 Btu/ft-hr-oF) of concrete
for the time constant calculation. The characteristic length of a
hollow cylinder is its wall thickness. The characteristic length
for the HI-STORM Overpack is therefore 29.5 in, or approximately
2.46 ft. Substituting into the equation, the overpack time constant
is determined as:
Error! Objects cannot be created from editing field codes.
One-tenth of this time constant is approximately 12.6 hours (756
minutes), substantially longer than the fire duration of 3.62
minutes, so the MPC is evaluated by considering the MPC canister as
an adiabatic boundary. The fuel temperature rise is computed next.
Table 4.III.130 lists lower-bound thermal inertia values for the
MPC-68M and the contained fuel assemblies. Applying design heat
load (36.942.8 kW (1.426x105 Btu/hr)) and adiabatic heating for the
3.62 minutes fire, the fuel temperature rise computes as:
F F hrTfuel °=
o
5
This is a very small increase in fuel temperature. Consequently,
the impact on the MPC internal helium pressure will be quite small.
Based on a conservative analysis of the HI-STORM 100 System
response to a hypothetical fire event, it is concluded that the
fire event does not adversely affect the temperature of the MPC or
contained fuel. We conclude that the ability of the HI-STORM 100
System to cool the spent nuclear fuel within design temperature
limits during and after fire is not compromised. An alternate
method using the FLUENT thermal model described in Section 4.III.4
can be adopted to evaluate HI-STORM site-specific fire accident
event similar to that described in Section 4.6 of HO-STORM FW
FSAR.
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(ii) HI-TRAC Fire1 To demonstrate the fuel cladding and MPC
pressure boundary integrity under an exposure to a hypothetical
short duration fire event during on-site handling operations, a
fire accident analysis of the loaded 100-ton HI-TRAC is performed.
This analysis, because of the lower mass of the 100-ton HI-TRAC,
bounds the effects for the 125-ton HI-TRAC. In this analysis, the
contents of the HI- TRAC are conservatively postulated to undergo a
transient heat-up as a lumped mass from the decay heat input and
heat input from the short duration fire. The rate of temperature
rise of the HI-TRAC depends on the thermal inertia of the cask, the
cask initial conditions, the spent nuclear fuel decay heat
generation, and the fire heat flux. Using conservatively bounding
inputs – lowerbound thermal inertia, steady state maximum cask
temperatures (Table 4.III.6) and design heat load (42.836.9 kW) - a
bounding cask temperature rise of 5.21178°F per minute is computed
from the combined radiant and forced convection fire and decay heat
inputs to the cask. During the handling of the HI-TRAC transfer
cask, the transporter is limited to a maximum of 50 gallons. The
duration of the 50-gallon fire using the methodology articulated
above for HI-STORM fire is 4.775 minutes. Therefore, the
temperature rise computed as the product of the rate of temperature
rise and the fire duration is 24.97°F, and the co-incident fuel
cladding temperature (734664.7oF)2 is below the 1058oF accident
limit. The elevated temperatures as a result of the fire accident
will cause the pressure in the water jacket to increase and cause
the overpressure relief valves to vent steam to the atmosphere.
Based on the fire heat input to the water jacket, less than 11% of
the water in the water jacket can be boiled off. However, it is
conservatively assumed, for dose calculations, that all the water
in the water jacket is lost. In the 125-ton HI-TRAC, which uses
Holtite in the lids for neutron shielding, the elevated fire
temperatures would cause the Holtite to exceed its design accident
temperature limits. It is conservatively assumed, for dose
calculations, that all the Holtite in the 125-ton HI-TRAC is lost.
Due to the increased temperatures the MPC experiences as a result
of the fire accident in the HI- TRAC transfer cask, the MPC
internal pressure increases. The pressure rise is computed using
the Ideal Gas Law and upperbound helium backfill pressure defined
in Chapter 4, Table 1.III.14.4.12 and results tabulated in Table
4.III.9. The computed MPC accident pressure is substantially below
the accident design pressure (Table 2.2.1). An alternate method
using the FLUENT thermal mode described in Section 4.III.5 can be
adopted to evaluate HI-TRAC site-specific fire accident event. (b)
Flood The flood accident is defined in Chapter 2 as a deep
submergence event. The worst flood from a thermal perspective is a
“smart flood” that just rises to the top of the inlets to prevent
airflow without
1 The HI-TRAC fire accident methodology is same as the generic
methodology in Section 4.6 of the HI-STORM 100 FSAR. 2 Computed by
adding the fire temperature rise to initial fuel temperature (Table
4.III.6).
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Proposed Rev. 143 REPORT HI-2002444 4.III-15
the benefit of MPC cooling by water. This effect is bounded by the
100% inlets ducts blockage accident evaluated herein in Section
4.III.6.2(d). (c) Burial Under Debris This accident event is
defined in Paragpragh 4.6.2.5. The methodology for tThe burial
under debris evaluation in Section 4.6 is employed to determine the
minimum available time for the fuel cladding to reach the accident
limit. Using the equation presented in Paragraph 4.6.2.5 and same
clad temperature margin presented in Table 4.6.6, burial time is
obtained and presented in Table 4.III.16. The coincident MPC
pressure bounding because of the following: (i) The MPC thermal
inertia is neglected. (ii) The initial storage temperatures under
MPC-68M storage are less than the HI-STORM 100 System
temperatures.is also computed and compared with the accident design
pressure (Table 2.2.1). The result (Table 4.III.16) is confirmed to
be below the permissible limit.
(d) 100% Blockage of Air Ducts This accident is defined in Section
4.6 as 100% blockage of the air inlet ducts for 32 hours. This
event is evaluated by blocking the air inlets in the FLUENT thermal
model and computing the 32- hour temperature rise of the MPC and
stored fuel. The results of this analysis are tabulated in Table
4.III.7. The results show that fuel cladding and component
temperatures remain below their respective accident limits
specified in Chapter 2 and Supplement 4.III. The increase in
temperature results in a concomitant rise of the MPC pressure. The
maximum accident pressure tabulated in Table 4.III.7 is below the
design limit specified in Chapter 2. Since the temperatures of
MPC-68M are bounded by the MPCs evaluated in Chapter 4, threshold
heat load defined in Table 4.6.8 can also be adopted for MPC-68M. A
threshold heat load is defined in Table 4.6.8 at or below which
periodic surveillance or vent blockage corrective actions are not
necessary. See section 4.6.2.4 for further details. (e) Extreme
Environmental Temperature The accident event is defined in
Paragraph 4.6.2.3. The principal effect of elevated ambient
temperature is a rise of the HI-STORM 100 temperatures from the
baseline normal storage temperatures by the difference between
elevated ambient and normal ambient temperatures. The results of
this event (maximum temperatures and pressures) are provided in
Table 4.III.17. The results are below the accident condition
temperature and pressure limits (Tables 2.2.1, 4.III.2 and 2.2.3).
As the normal storage temperatures under MPC-68M storage in the
HI-STORM 100 overpack are bounded by the HI-STORM 100 System
temperatures reported in Section 4.4, the temperatures under this
event are likewise bounded by the extreme ambient evaluation in
Section 4.6. (f) 100% Rods Rupture Accident
ATTACHMENT 5 TO HOLTEC LETTER 5014811
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HOLTEC INTERNATIONAL COPYRIGHTED MATERIAL HI-STORM 100 FSAR
Proposed Rev. 143 REPORT HI-2002444 4.III-16
In accordance with NUREG-1536 a 100% rods rupture accident is
evaluated assuming 100% of the rods fill gases and fission gases
release in accordance with NUREG-1536 release fractions. The
MPC-68M pressure under this postulated accident is computed and
tabulated in Table 4.III.4. The pressure is below the accident
design pressure (Table 2.2.1). (g) Jacket Water Loss The principal
effect of jacket water loss accident is a temperature increment in
the stored fuel and MPC from the baseline conditions under in a
HI-TRAC. As the MPC-68M temperatures in the HI- TRAC are bounded by
MPC-68 32 temperatures (see Table 4.5.6) the jacket water loss
temperatures are likewise bounded by the HI-TRAC jacket water loss
evaluation in Section 4.6. 4.III.7 REGULATORY COMPLIANCE As
required by ISG-11, the fuel cladding temperature at the beginning
of dry cask storage is maintained below the anticipated
damage-threshold temperatures for normal conditions for the
licensed life of the HI-STORM System. As required by NUREG-1536
(4.0,IV,3), the maximum internal pressure of the cask remains
within its design pressure for normal, off-normal, and accident
conditions. Design pressures are specified in Table